THE primary function of a thermal protection system (TPS) is to
|
|
- Jocelyn Gilbert
- 5 years ago
- Views:
Transcription
1 AIAA JOURNAL Vol., No., Feruary 1 Two-Dimensional Orthotropic Plate Analysis for an Integral Thermal Protection System Oscar Martinez, Bhavani Sankar, and Raphael Haftka University of Florida, Gainesville, Florida 311 and Max L. Blosser NASA Langley Research Center, Hampton, Virginia 381 DOI: 1.1/1.J11 This paper is concerned with homogenization of a corrugated-core sandwich panel, which is a candidate structure for integrated thermal protection systems for space vehicles. The focus is on determining the local stresses in an integrated thermal protection system panel sujected to mechanical and thermal loads. A micromechanical method is developed to homogenize the sandwich panel as an equivalent orthotropic plate. Mechanical and thermal loads are applied to the equivalent thick plate, and the resulting plate deformations were otained through a shear deformale plate theory. The two-dimensional plate deformations are used to otain local integrated thermal protection system stresses through reverse homogenization. In addition, simple eam models are used to otain local facesheet deformations and stress. The local stresses and deflections computed using the analytical method were compared with those from a detailed finite element analysis of the integrated thermal protection system. For the integrated thermal protection system examples considered in this paper, the maximum error in stresses and deflections is less than %. This was true for oth mechanical and thermal loads acting on the integrated thermal protection system. Nomenclature a = integrated thermal protection system length = integrated thermal protection system width d = height of sandwich panel (centerline to centerline) D = inverse ending stiffness matrix D = reduced ending stiffness matrix E = Young s modulus EI = equivalent flexural rigidity G = transverse shear modulus P z = pressure load acting on the two-dimensional orthotropic panel fqg = displacement vector Q ij = transformed lamina stiffness matrix Q x, Q y = shear force on unit cell Q ij = transformed lamina stiffness matrix in the nonprincipal coordinate system s = we length t BF = ottom facesheet thickness t TF = top facesheet thickness t w = we thickness T D e = deformation transformation matrix of the ith component of the corrugated core U = unit-cell strain energy v = deflection of eam w = integrated thermal protection system panel deflection Received 13 January 11; revision received 9 August 11; accepted for pulication 31 August 11. Copyright 11 y the American Institute of Aeronautics and Astronautics, Inc. Copies of this paper may e made for personal or internal use, on condition that the copier pay the $1. per-copy fee to the Copyright Clearance Center, Inc., Rosewood Drive, Danvers, MA 193; include the code 1-1/1 and $1. in correspondence with the CCC. Structural Analyst, Department of Mechanical and Aerospace Engineering; Jacos Engineering, Engineering Science Contract Group, Houston, Texas. Eaugh Professor, Department of Mechanical and Aerospace Engineering. Associate Fellow AIAA. Distinguished Professor, Department of Mechanical and Aerospace Engineering. Fellow AIAA. Aerospace Technologist, Metals and Thermal Structures Branch, Structural Concepts and Mechanics Branch. 38 y f = local y axis of facesheet y w = local y axis of we z f = local z axis of facesheet z w = local z axis of we " o = midplane strain = angle of we inclination = curvature = roots of auxiliary equation = Poisson s ratio x, y = rotations of the plate s cross section p = unit-cell length I. Introduction THE primary function of a thermal protection system (TPS) is to protect the space vehicle from extreme aerodynamic heating and to maintain the underlying structure within acceptale temperature and mechanical constraints [1]. The current state-of-the-art reusale TPS covers the outer surface of a vehicle with a layer that does not carry significant loads ut prevents the underlying loadearing structure from exceeding acceptale temperature limits. An intriguing alternative is to emed the relatively fragile insulation in a structural load-earing sandwich panel. A roust, structural facesheet would then form the outer surface of the vehicle, therey potentially removing some vehicle operational constraints, e.g., flight through rain, and reducing required maintenance. By using the insulation thickness as a structural sandwich core, there is a potential to increase structural ending stiffness, save weight, and reduce overall thickness of the vehicle wall. There are significant challenges in designing a structure to function simultaneously as an efficient aerospace load-earing structure and a thermal insulator. The challenges include understanding the key thermal-structural loads and requirements that drive the design, identifying materials that est perform in a thermal environment, and accommodating thermal expansion mismatches, thermal stresses, and effects of temperature-dependent properties inherent in these systems. Candidate concepts comprising metallic foams and innovative core materials, such as corrugated and truss cores, have een investigated. The corrugated-core sandwich panel is composed of several unit cells placed adjacent to each other. The empty space in the corrugated core will e filled with a non-load-earing insulation such as Safill. The TPS concept comines all three passive TPS concepts: heat sink,
2 388 MARTINEZ ET AL. hot structure, and insulated structure []. The sandwich structure will replace the structural skin and the insulation in the current thermal structures. Since the sandwich construction is stiffer than single skin construction, the numer of frames and stringers will e significantly reduced. Furthermore, the insulation is protected from foreign oject impact and requires less or no maintenance such as waterproofing. The integral TPS/structure [integrated thermal protection system (ITPS)] design (Fig. 1) can significantly reduce the overall weight of the vehicle as the TPS/structure performs the load-earing function. The ITPS is expected to e multifunctional (offer insulation as well as load-earing capaility). Advantages of an ITPS concept have een identified y Martinez et al. [3]. The finite element method () is commonly used to perform three-dimensional (3-D) stress analysis of the ITPS structure. However, a full-scale 3-D finite element (FE) stress analysis will e expensive and time-consuming for a quick preliminary sizing analysis of an ITPS. The same is true with design optimization, which may involve 1s of analyses. An alternative is to model the ITPS as an equivalent orthotropic plate and perform a -D plate analysis to otain the displacement and stress fields. This paper is the third in the sequel of papers written y the authors on the suject of homogenization of the ITPS. The procedures are depicted in Fig. 1a. The ITPS, which consists of repeating cells, is homogenized as an equivalent orthotropic plate, and the ABD matrices (stiffness matrices) of the equivalent plate are calculated using analytical methods. In the first paper [3], the authors descried the homogenization procedures and verified the accuracy of the extensional A, ending D, and shear C stiffness matrices y comparing with those calculated using s. In the second paper [], procedures for calculating equivalent thermal forces and moments for a given through-the-thickness temperature distriution were developed without considering shear deformations. In the current paper, the equivalent orthotropic plate is analyzed using a first-order shear deformale theory in which the transverse shear stresses are not neglected. Both pressure loads and thermal loads are considered. The plate analysis results include midplane strains f" g and curvatures fg at a given x; y. Then, a reverse homogenization procedure is used to calculate the detailed stress field from the f" g and fg otained from the plate analysis. The stresses are compared with those otained from a detailed 3-D analysis of the ITPS. It is worth pointing out that, in the present work, analytical methods are used for every step of the procedures descried aove. methods are much faster than FE analysis and are suitale when 1s of analyses have to e performed in the design optimization of the ITPS panel. methods will e extremely useful if one decides to use stochastic optimization rather than a safety-factor-ased design approach. Only a limited amount of work in micromechanical unit-cell analyses of sandwich structures is availale in the literature [ 1]; it has een summarized in our previous works [3,] and will not e repeated here. The aforementioned researchers used the forcedistortion relationship, which involved mechanics of materials equations to determine average unit-cell stresses. In the current paper, we improve on that y use of a first-order shear deformale plate theory (FSDT) [13] for the -D equivalent ITPS plate analysis, reverse homogenization, and simple eam models for the local ITPS stress analysis. The higher-order theory that includes the effects of transverse shear deformation is suitale for an ITPS ecause of the significant transverse shear loads that are carried y the corrugated core. Although such stresses can e postcomputed through 3-D elasticity equilirium equations, they are not always accurate. The classical laminate plate theory (CLPT) [1] underpredicts deflections and overpredicts uckling loads with plate length-to-thickness ratios less than. For this reason alone, it was necessary to use the Fig. 1 ITPS: a) corrugated-core sandwich panels for use as an ITPS and ) flowchart depicting procedures in homogenization of the ITPS panel and reverse homogenization to otain the detailed displacements and stresses (TFS denotes top facesheet; BFS denotes ottom facesheet).
3 MARTINEZ ET AL. 389 Fig. Dimensions of the unit cell. the in-plane forces, transverse shear forces, and ending and twisting moments; and A, C, and D are the extensional, transverse shear, and ending stiffness. The equivalent orthotropic plate is assumed to e symmetric aout its midplane, leading to a negligile extensionending coupling matrix B. The case of unsymmetric plates will e dealt with using approximate methods as discussed in this paper. The detailed homogenization procedures for determining the A, B, C, and D matrices for the corrugated-core sandwich panel were derived in Martinez et al. [1]. In the following sections, we descrie the procedures for determining the response of the ITPS to pressure and thermal loads. first-order theory in the ITPS analysis of relatively thick plates. The primary ojective and importance of using the FSDTwhen analyzing the ITPS as a two-dimensional (-D) homogenized plate was to ring out the effect of shear deformation on deflections and stresses. The paper s ojective is to derive a satisfactory approximate plate theory of an ITPS panel homogenized as a -D plate using the FSDT. The solution of the plate equations gives all necessary information for calculating stresses at any point of the thick plate. Local ITPS stresses of the facesheets and wes are derived y using the -D plate deflections w, rotations, x, y, reverse homogenization, and a transformation matrix. Reverse homogenization only captures the midplane strains and curvatures of the unit cell at the origin (see Fig. ), which resulted in a uniform stress field in the facesheets and wes. Additional local analysis had to e done through one-dimensional (1- D) eam models to capture the nonuniform stress field of the facesheet and wes when sujected to mechanical and thermal loads. The ITPS panel is sujected to a uniform pressure load and thermal edge moments, and the resulting plate deflections and local ITPS stresses are compared with a detailed FE analysis of the ITPS. II. Analysis Consider a simplified geometry of an ITPS unit cell, shown in Fig.. The z axis is in the thickness direction of the ITPS panel. The stiffer longitudinal direction is parallel to the x axis, and the y axis is in the transverse direction. The unit cell consists of two inclined wes and two thin facesheets. The unit cell is symmetric with respect to the yz plane. The upper faceplate thickness t TF can e different from the lower faceplate thickness t TB, as well as the we thickness t w. The unit cell can e identified y six geometric parameters p; d; t TF ;t BF ;t w ; (Fig. ). The equivalent stiffness of the orthotropic plate is otained y comparing the ehavior of a unit cell of the corrugated-core sandwich panel with that of an element of the idealized homogeneous orthotropic plate (Fig. 3). The in-plane extensional and shear response and out-of-plane (transverse) shear response of the orthotropic panel are governed y the following constitutive relations: 3 3( ) N A 3 3 "o Q 3 C 3 or ffgkfqg (1) M 3 3 D In Eq. (1), " o and are the in-plane and transverse shear strains, and are the ending and twisting curvatures; fng, fqg, and fmg, are A. Uniform Pressure Loading A simply supported orthotropic sandwich panel of width (y direction) and length a (x direction) was considered for a -D plate analysis (Fig. 3). The simply supported oundary conditions are descried as w;y; wa; y M x ;y M x a; y wx; ; wx; M y x; ; M y x; () The panel is exposed to a uniform pressure load that can e represented y the doule Fourier sine series, as shown in Eq. (3): P Z x; y X1 X 1 mx ny P mn sin sin (3) a m1 n1 In the aove equation, P mn 1P o = mn for uniform pressure loads [13], where P o is the intensity of the uniformly distriuted load. The ITPS panel is also assumed to have the following deformations that satisfy the simply supported oundary conditions: wx; y X1 X 1 mx ny A mn sin sin (a) a xx; y X1 yx; y X1 m1 n1 X 1 m1 n1 X 1 m1 n1 mx B mn cos a sin ny mx ny C mn sin cos a () (c) In the aove equations, wx; y is the out-of-plane displacement, and x x; y and y x; y are the plate rotations. The unknown constants A mn, B mn, and C mn were otained y sustituting the constitutive relations in the form of the assumed deformations from Appendix A [Eqs. (A) and (A3)] into the differential equation of equilirium [Eq. (A1)]. This results in a system of three linear equations for the unknown coefficients as shown: L P o A I B II C L 1 III D z y Fig. 3 Equivalent orthotropic thick plate for the unit cell of the corrugated-core sandwich panel. Fig. Half of the top face under the action of a uniform pressure loading.
4 39 MARTINEZ ET AL. m D 11 a n D mn kc D 1 D a mn n D 1 D D a m D a kc kc kc m a kc n m kc a n kc m a kc n 3 ( Bmn C mn A mn ) ( ) P mn () After solving for the unknown constants, the deflections were otained at a given x; y coordinate on the -D orthotropic sandwich panel using Eq. (). The results in the aove series converged for m n 3, resulting in a total of 9 terms in Eq. () [1]. It should e mentioned that the shear correction factor k in Eq. () was assumed to e unity, as C and C were directly predicted using micromechanics [3]. B. Top Facesheet Local Deflection The top facesheet was directly exposed to a uniform pressure loading. The uniform pressure loading that acted on the thin facesheet resulted in local deflections that were not captured y the -D FSDT. The FSDT analysis resulted in midplane strains and curvatures at the center of the unit cell. The facesheet local deformation vector due to the pressure requires additional analysis. A 1-D eam prolem was used to calculate the additional local deformation of the facesheet under uniform pressure load. Because of symmetry, only half of the top facesheet was analyzed (Fig. ). A matrix structural analysis method (FE) was used to determine the deflection of the top facesheet. Half of the top facesheet was modeled using two eam elements. Displacement and rotation oundary conditions were imposed on the two ends of the facesheet. A torsional spring element with a rotation oundary condition was used at the middle node to capture the resistance to rotation offered y the we connected to the facesheet. The elements are indicated y roman numerals in Fig.. Because of symmetry, A and c were zero. Because of the displacement constraint of the we and the facesheet, the deflection at junction B was set to zero, i.e., v B. The clamped oundary condition at node D resulted in D eing zero. The consistent, or work, equivalent loads on each element due to the uniform pressure loading were computed using the eam element shape functions [1] as shown elow: 8 >< >: F i C i F j C j 9 >= >; Z L P o 8 >< y f y f >: 1 3y f y3 f L L 3 y3 f L L 3y f y3 f L L 3 y f L y3 f L 9 >= dy f () In Eq. (), i and j correspond to the left and right nodes of the eam and torsional elements, and L corresponds to the length of the eam element. The element stiffness matrix for the eam and torsional spring elements were assemled accordingly to otain the gloal stiffness matrix and gloal force vector: EI 1 L L 3 L 1 8 >< >: L 1 L P ol 1 P ol P o L 1 P ol 1 1 L 1 L 1 1 L 1 L 3 9 >= >; 3 8 >< >: v A v C B 9 >= >; >; KfqgfFg () In Eq. (), L 3 was the length of the we. Equation () was solved to otain the nodal displacements, which were sustituted ack in the interpolation functions to otain the transverse deflection of the facesheet: v I y f 1 3y f L y3 f v L 3 A y f L y3 f L B (8) v II y f y y f L y3 f 3y f L B L y3 f v L 3 C (9) The 1-D deflections from Eqs. (8) and (9) were superimposed with the plate deflection from Eq. (a). C. Stresses due to Pressure Loading 1. Stress Because of Midplane Strain and Curvature The FSDT analysis of the equivalent plate yields midplane strains, curvatures, and transverse shear strains at given x; y of the equivalent plate. These deformations are called the macrodeformations [3], and they are the deformations in the equivalent orthotropic plate. To calculate the stresses in the ITPS, we apply the aforementioned macrodeformations to a unit cell situated at the same point. Then, using the transformation matrices defined y Martinez et al. [3], the (actual) deformations in the facesheets and the we were calculated. From the constitutive relations of the facesheet and we materials (layup in the cases of laminates), the stresses are calculated. A flowchart (Fig. ) outlines the stress calculation procedure. In the flowchart, the x and y locations correspond to the ITPS -D equivalent plate, and y and z correspond the local location on either the face or we; Fig. a. If e in Fig. was equal to 3 or (left or right we), then the micro-mid-plane strains and curvatures were a Fig. Local stress flowchart for an ITPS as a -D plate with transverse shear force effects consideration.
5 MARTINEZ ET AL. 391 Fig. Unit cells: a) ITPS unit cell under the action of transverse shear force, and ) half-itps unit cell. function of y w, which would result in we stresses as a function of y w. There will e additional stresses in the top facesheet that result from the local effects of the pressure loading. These local stresses are determined from the local deflections derived in Eqs. (8) and (9) and added to that resulting from the equivalent plate analysis in Eq. (). The stresses calculated from Eqs. (8) and (9) neglect a shear deformation term ecause the facesheets and wes are thin when considered y themselves and compared with the unit-cell dimensions.. Stress due to Transverse Shear Strains In this section, we descrie the procedures to calculate the stresses in the facesheets and we due to transverse shear force Q y acting on the equivalent orthotropic plate. Figure illustrates the ITPS halfunit cell under the action of various forces that resulted from a transverse shearing force Q y. The unknown forces P, R, and F for unit Q y were determined from Castigliano s second theorem [1]. Then, using the procedure descried y Martinez et al. [1], the ending moments at a cross section of the facesheets or wes were determined (Appendix B). From the moment equations (Appendix B), the resulting curvature in either of the facesheets or wes was determined: eq y D e 1 y y; x; y D e 11 D e M ij yq y x; y (1) The flowchart shown in Fig. takes into account the moments from Eq. (A1) and the transverse shear force from Eq. (A3) for local stress determination. The superscript e in Eq. (1) refers to a component assignment of the ITPS. An e of 1,, 3, and refers to the top facesheet, ottom facesheet, left we, and right we. It should e mentioned that, in applying Castigliano s theorem, we accounted only for the strain energy due to ending in various segments of the cross sections. The energy due to in-plane forces and transverse shear force was assumed to e negligile compared with the ending energy since the thickness of the facesheet and wes are small when compared with its length. 3. Local Stress due to the Uniform Pressure Loading: One-Dimensional Analysis The calculation of local stresses in the top facesheet due to pressure loading is a continuation of the analysis descried aove that determines the local transverse displacements [see Eqs. (8) and (9)]. The free ody diagram of the top facesheet under the action of a uniform pressure loading is shown in Fig.. In Fig. 8, the moments at points A and B were determined y multiplying the element stiffness matrix with the corresponding node displacement vector; see Appendix C. The stresses at any local point of the top facesheet were determined y multiplying the moment vector with the inverse of the ending stiffness matrix and the lamina stiffness matrix of the facesheet as shown elow. A superscript value of 1 refers to the top facesheet: 3 xx y f ; z f yy y f ; z f 1 yy y f ; z f z f Q 1 11 Q 1 1 Q 1 1 Q 1 1 Q 1 Q 1 Q 1 1 Q 1 Q 1 3 D 1 11 D 1 1 D 1 D 1 1 D 1 1 D 1 D 1 1 D 1 D D 1 1 D 1 M ij y M ij y 3 (11) D. Unsymmetric Integrated Thermal Protection System Panel Configuration The methods descried so far are for an ITPS with symmetric facesheet configurations with respect to the unit-cell y axis, i.e., B matrix is zero. However, certain designs may call for different materials or thicknesses for the top and ottom facesheets that make it perform at its optimum for mechanical and thermal applications. Ceramic matrix composites and high-temperature metals are possile choices for the top facesheet due to their high service temperatures. Aluminum alloys and composite materials are of preference for the ottom facesheet due to their high specific heats and mass-efficient structural properties. An ITPS with different materials for the top and ottom facesheets results in a stiffness matrix with a nonzero coupling stiffness matrix, B. This type of ITPS configuration will e referred to as an unsymmetric ITPS. The presence of the ending-extensional coupling stiffness matrix causes considerale difficulty in otaining solutions to a plate prolem using either the FSDT or CLPT method. As an alternative, the reduced stiffness matrix approach [1], wherein B is assumed to e equal to zero and D is modified, could e used. Then, the techniques descried aove for analysis of orthotropic plates can e directly used. The reduced stiffness matrix [Eq. (1)] was used in the plate solution of an orthotropic plate with a uniform pressure load. The use of the reduced stiffness matrix tends to reduce the effective stiffness P o y f y f P L + y ) o( f ( L + y f ) M C M AB C M C B C M BC L y f y f F B Fig. Free ody diagram of section BC (see Fig. ) of the top facesheet. Fig. 8 Free ody diagram of section AB (see Fig. ) of the top facesheet.
6 39 MARTINEZ ET AL. of the plate, which results in increased ending deflections while uckling loads are decreased when compared with equivalent symmetric orthotropic plates. The corresponding error in the out-ofplane plate displacement and local stresses was investigated and compared with the FE results. The reduced ending stiffness matrix D is given y [1]: D DBA 1 B (1) E. Thermal Stress Moment Resultants The ITPS panel is exposed to extreme reentry temperatures that result in various temperature distriutions through the ITPS thickness as a function of reentry time (transient heat transfer) [1]. The temperature distriution causes the ITPS to deform due to the resulting thermal force and moments that were determined y Martinez et. al [18]. An analytical solution was derived to predict the thermal deflection of the ITPS panel when sujected to a throughthe-thickness temperature distriution. A -D plate analysis was needed to determine the response of the ITPS when it was sujected to thermal moments. Consider a rectangular plate simply supported along the nonloaded edges and ent y moments distriuted along the edges x ;a (Fig. 9). The distriuted moment in Fig. 9 is equal to the previously calculated thermal moment from Martinez et al. [18]. Typical ranges of L=h for the ITPS will e from 1 to, which result in nonnegligile shear effects. An FSDTapproach was implemented for the deflection solution of the ITPS plate. The oundary conditions for the thermal plate prolem are as follows: w ; x y ; w ; y x ;a Mx T D 11 x;x x ;a (13) The deflections and rotations are assumed such that they satisfy all displacement oundary conditions along the edges y and y : We assume solutions to e of the form of an exponential series function; see Eq. (D). Sustituting the assumed solution from Eq. (D) in the differential equations [Eq. (D1)] yields a set of homogeneous linear equations for the unknown coefficients,, and as shown elow: D 11 D A D 1 D A 3 det D 1 D D D A A A A A A 89 < = 89 < = : ; : ; (1) The lamdas were solved y setting the determinant of the coefficient matrix to zero. The solution to Eq. (1) yielded six solutions ;. The six roots yielded six terms in Eq. (D), which resulted in three equations with six unknowns per equation [Eqs. (D3 D)]. Eighteen unknown constants resulted from the equations in Appendix D, which were not possile to solve with only six oundary conditions. However, a relation was made etween all 18 constants (see Appendix D), which reduced the numer of independent constants from 18 to. A unique solution to each unknown constant was now availale ecause of the equal numer of unknown constants and oundary conditions. The moment distriution along the edges of the plate was represented y the Fourier sine series as M x XN n1 M n sin nx a (1) For the case of a uniform distriution of the ending moments, the term M n was represented as M n M T x =n. A system of six linear equations was otained y sustituting Eq. (D) into Eq. (1) and Eq. (1) into Eq. (13): 8 >< >: >= >; e 1 e e 3 e e e C 1 C C 3 C C C C 1 e 1 C e C 3 e 3 C e C e C e D D 1 1 D D D D 1 3 D D 1 D D 1 D D 1 D D 1 1 e 1 D D 1 1 e 1 D D 1 1 e 1 D D 1 1 e 1 D D 1 1 e 1 D D 1 1 e >< >: M T y m My T m 9 >= >; (1) wx; y XN n1 A n x sin ny yx; y XN n1 xx; y XN C n x cos ny n1 B n x sin ny (1) By solving the set of six linear equations, the assumed deflection equations [Eq. (1)] were solved. Equations (D) and (D) were used to otain the solutions to the plate rotation in the x and y directions. The same procedure was applicale for uniformly distriuted The deflection and curvature solutions of the plate are taken in the form of a series where A n x, B n x, and C n x are all unknown functions of x only. It was assumed that the two unloaded edges are simply supported; hence, each term of the series satisfies the oundary conditions w and x on these two sides. It remains to determine A n x, B n x, and C n x in such a form to satisfy the oundary conditions on the sides x and x a, and the equation of equilirium [Eq. (A1)]. Taking the assumed deflection and rotations in the form of a series and sustituting them into the equilirium equations results in three ordinary differential equations [Eq. (D1)]. h x L T M x Fig. 9 ITPS orthotropic plate sujected to uniformly distriuted thermal end moments. y a
7 393 MARTINEZ ET AL. Fig. 1 ITPS out-of-plane deformation. FE a) model and ) mesh. III. Results A. Integrated Thermal Protection System Out-of-Plane Displacement: Pressure Load For verification of the accuracy in prediction of the deflection and stress results of an ITPS sandwich panel under the action of a uniform pressure loading, an ITPS panel with the following dimensions was analyzed: p mm, d 1 mm, ttf 1: mm, tbf 1: mm, tw 1: mm,, a 1 m, and 1 m. An Inconel 1 alloy was used as an example to verify the analytical models for determining local stresses in the ITPS sandwich panel (E GPa, :8). A 1-unit-cell ITPS sandwich panel was modeled for the analysis (Fig. 1a). Because of symmetry of the ITPS panel, only a quarter of the plate was modeled. A 3-D FE analysis was conducted on the ITPS plate using the commercial ABAQUSTM version. FE program. Eight node shell elements were used to model the facesheets and wes of the unit cell. The shell elements have the capaility to include multiple layers of different material properties and thicknesses that are needed to model a laminate. Three integration points were used through the thickness of the shell elements. The model consisted of 1,8 nodes and 1 elements (Fig. 1). The ITPS panel was sujected to a uniform pressure loading of 8,9 Pa (1 psi), and symmetric oundary conditions were imposed at the edges of the ITPS plate. The ottom facesheet had an out-of-plane displacement constraint w ; y w x;, while the top facesheet only had a rotation constraint in all three directions along x and y (Fig. 11). The FE model was constrained with the appropriate oundary conditions from Eq. () and appropriate symmetric oundary conditions. The ysymmetric oundary conditions are v x; = Rx x; = Rz x; = and the x-symmetric oundary conditions are potentially have an unlimited numer of fictitious unit cells depending on what x and y points are chosen. Since the ITPS model contained 1 unit cells, for comparison, the homogenized ITPS -D plate was divided into a 1 1 grid. The accuracy was verified through FE results and a convergence study. Each intersection of the grid lines corresponded to an analytical ITPS unit cell that corresponded to the center location of the unit cell. The curvatures in the x and y directions at each analytical grid intersection were assumed to e constant for the entire unit cell; however, the curvatures in the x and y directions from the 3-D results were not constant within each unit cell. The curvatures varied through each unit cell. This stress gradient effect could not e captured y the analytical homogenized plate model. The out-of-plane displacement contour of the ITPS is illustrated in Fig. 1. The out-of-plane displacement results at x a= were compared with Out-of-Plane Displacement (mm/po) ending moments along the y axis. In the case of simultaneous action of couples along the entire oundary of the ITPS plate, the deflections and moments can e otained y suitale superposition of the results otained from the detailed discussion. x 1-8 FSDT CLPT Length Along x=a/ (m) Out-of-Plane Displacement (mm/po) Fig. 1 x 1-9 FSDT CLPT Length Along x=a/ (m) Fig. 13 Out-of-plane displacement comparison etween and analytical solution of the top and ottom facesheets. Out-of-Plane Displacement (mm/po) The out-of-plane displacements of the ITPS panel at x a= were extracted and compared with the results otained from Eqs. (a), (8), and (9). The homogenized ITPS -D plate can Out-of-Plane Displacement (mm/po) u a=; Rx a=; Rz a=; Fig. 11 FE oundary conditions for the ITPS plate. x Length Along x = a / (m) x Length Along x = a / (m) Fig. 1 Out-of-plane displacement comparison etween and analytical solution of the top and ottom facesheets for an unsymmetric ITPS.
8 39 MARTINEZ ET AL. the analytical results (Fig. 13). As seen in Fig. 13, the analytical outof-plane displacement results are within a 1 % difference from the results. There was greater percentage difference near the oundary of the ITPS ecause the oundary conditions created localized effects that were not captured y the FSDT method and the 1-D eam analysis. At the center of the ITPS plate, the structure acted more like the homogenized -D plate ecause it was farther away from the oundary. The superposition of the displacement results otained from the FSDT method and the 1-D eam analysis resulted in a less than % difference and an accurate representation of the top facesheet deflection response when sujected to a uniform pressure load. The ottom facesheet FE displacement results agreed well with the displacement results otained from the FSDT method. The ottom facesheet deflection acted more like a -D plate ecause of the asence of local effects from the uniform pressure loading. To verify the applicaility of present methods to unsymmetric ITPS, an ITPS with a different ottom facesheet was considered. The ottom facesheet was assumed to e of an aluminum alloy (E :19 GPa, 1 :3). An ITPS panel with the same dimensions as mentioned aove was modeled and analyzed for investigation of the error in using the reduced stiffness matrix in the FSDT method for an unsymmetric ITPS panel. The local stresses were extracted at x a= and compared with the analytical stresses that resulted from the FSDT method y use of the reduced stiffness matrix. From Fig. 1, it can e noted that the plate displacement results were not affected y using the reduced stiffness matrix in the FSDT plate analysis. The largest difference etween the and analytical results was less than %. Normalized Core Thickness 1 Tale 1 Thermal moments of Inconel 1 under the s reentry temperature distriution Stiffness M x, Nm=m M y, Nm=m 1:E 8:8E FE 1:1E 9:E % diff..1. s -1 8 Temperature (K) Fig. 1 ITPS temperature distriution at s reentry time. Boundary condition 1 Boundary condition Boundary condition 3 Tale Boundary conditions of the ITPS panel Top face rotation Top face displacement Bottom face rotation Bottom face displacement X X X X B. Integrated Thermal Protection System Out-of-Plane Displacement: Temperature Distriution From the results otained y Martinez [19] on the analysis of a plate under the action of thermal edge moments, confidence was gained with the analytical procedure for a plate solution of an orthotropic thick plate under the action of uniformly distriuted edge moments. The next step was to apply that approach to analytically predict the out-of-plane displacement of an ITPS under the action of thermal moments from an applied temperature distriution. The temperature distriution that was considered for the analysis was at the s reentry time of a space-shuttle-type vehicle, as descried in [19] (Fig. 1). The ITPS was composed of Inconel 1, and the resulting thermal moments for Inconel 1 are calculated and compared with the FE verification procedure discussed y Martinez [19]. The greatest percentage difference etween the analytical and comparison was less than % for M T x. The thermal moment results (Tale 1) were used to determine the analytical out-of-plane displacement of the ITPS panel. Tale 1 consists of two moments. The aforementioned thermal plate analysis with edge moments on the x and y axes can e used through suitale superposition (see Fig. 1). The same ITPS panel that was modeled for the pressure load example was used for the temperature distriution example, except for this case, the thickness of the faces and wes was increased from 1. mm (.9 in.) to. mm (.8 in.). The numer of nodes and elements remained the same, as well as the oundary conditions. The s temperature distriution was included in the input file, and the out-of-plane displacement of the ITPS was extracted at x a= after analysis. There was no need to do any local out-ofplane displacement prolem for this case ecause the temperatures of the top and ottom facesheets were considered to e constant ecause of the thin faces when compared with the ITPS thickness. Therefore, the faces were only under the action of a thermal expansion. The out-of-plane displacement result of the ITPS panel under the action of the fourth-order temperature distriution is shown in Fig. 1 along with the CLPT [19] and FSDTanalytical results. The FSDT method for predicting out-of-plane displacement for an ITPS under the action of uniformly distriuted edge moments does not predict accurate results for this particular ITPS panel. The FSDT method overpredicts the actual FE displacement y %, and the CLPT [] method underpredicts the actual FE displacement y X X X ITPS Out-of-Plane Displacement (mm) 8 CLPT FSDT TFS- BFS- Fig Length Along x = a / (m) a) ) ITPS a) out-of-plane thermal displacement and ) thermal displacement contour (TFS denotes top facesheet; BFS denotes ottom facesheet).
9 MARTINEZ ET AL. 39 ITPS Out-of-Plane Displacement (mm) 3 1 TF1 CLPT BF1 TF BF TF3 BF3 FSDT Length Along x = a / (m) Fig. 1 ITPS out-of-plane displacement due to a fourth-order temperature distriution for various edge oundary conditions. ITPS Center Panel Displacement (mm) 1 1 FSDT % difference 1 1 Length / Height (L / h) Fig. 18 ITPS center panel displacement for various L=h values. %. The edge oundary conditions of the top and ottom facesheets of the model were modified in hopes of otaining a good comparison etween the and analytical results. The set of oundary conditions are listed in Tale. An X represents a constrained condition. The ITPS out-of-plane displacement results for all different oundary conditions are shown in Fig. 1 and compared with the CLPT and FSDT results. As seen from Fig. 1, the oundary condition does not affect the center plate out-of-plane displacement. The different oundary conditions had an effect on the out-of-plane displacements that were near the edge oundary. The different oundary conditions cause different oundary effects that affect the ITPS out-of-plane displacement that cannot e captured y the FSDT or CLPT methods. The change in oundary condition did not lead to a etter comparison etween the and analytical results Percentage Difference % σ xx (MPa/Po) σ yy (MPa/Po) x Length at x = a / (m) x Length at x = a / (m) Fig. Top facesheet stress in the x and y directions of the ITPS panel. The thermal moments that resulted from the fourth-order temperature distriution caused the ITPS panel to end due to thermal stresses. Since the ITPS plate had a low L=h value, the thermal moments acted on a short edge length of the ITPS, which introduced other local oundary effects into the results. These local ITPS plate effects that resulted from the thermal moments were not fully captured y the FSDT and CLPT methods when the plate had a low L=h ratio. The L=h ratio of the ITPS was increased to investigate what L=h ratio of an ITPS will result in a less than a % prediction error of the and analytical out-of-plane thermal displacement results. The L=h ratio in the model was increased y adding more unit cells to the ITPS plate while keeping the ITPS thickness constant. For ITPS plates with large L=h ratios, the thermal moments now acted on a long edge length, and all the local ITPS oundary effects were minimized. The ITPS panel was ale to end as a plate, and the ending ehavior was captured y the FSDT that was outlined aove. The resulting maximum out-of-plane displacement was extracted from the output after analysis and compared with the maximum analytical plate displacement for various L=h ratios. The percentage error decreased from % for low L=h ratios to % for high L=h ratios (Fig. 18). The analytical and results were in good agreement for that particular L=h ratio, which resulted in a less than % prediction error of the ITPS out-of-plane displacement from a fourth-order temperature distriution (see Fig. 19). The and analytical stress results were plotted at x a= for an ITPS panel with an L=h 18 (Figs. and 1). C. Integrated Thermal Protection System Local Stress The local stresses in the x and y directions were extracted from the output after analysis at x a=. The analytical top facesheet stresses were computed y superimposing the stresses results from Fig. and Eq. (1). The ottom facesheet and we stresses were computed from Fig.. All stress results were computed at z f or w t=, which was the ottom surface of each component. Similar to the out-of-plane displacement results, the percentage difference etween the and analytical results was greater at the oundary of the ITPS; see Figs.. The percentage difference near the oundary of the ITPS was 9%. The percentage difference ITPS Out-of-Plane Displacement (mm) 1 BF- TF- FSDT....8 Length Along x = a / (m) Fig. 19 ITPS out-of-plane displacement due to a temperature distriution for an ITPS with L=h 18. σ xx (MPa/Po) σ yy (MPa/Po) 1 x Length Along x = a / (m) x Length Along x = a / (m) Fig. 1 Bottom facesheet stress in the x and y directions of the ITPS panel.
10 39 MARTINEZ ET AL. Fig. Top facesheet stress in the x and y directions of an unsymmetric ITPS panel. σ xx (MPa/Po) σ yy (MPa/Po) x Length Along x = a / (m) 1 x Length Along x = a / (m) Fig. 3 Bottom facesheet stress in the x and y directions of an unsymmetric ITPS panel. was less than % for stress results near the center of the ITPS plate. The percentage difference for the we stress in the x direction was less than 1% when compared with the results. The percentage difference for the we stresses in the y direction was less than 3% at the center of the plate and greater than % near the oundary of the ITPS panel. The analytical local stress of the right we in the y w direction was different than the results ecause, near the oundary of the ITPS, there was a large change in curvature due to the local oundary effect that was not captured y the analytical model. The curvature was changing through the length of the unit cell while, in the analytical model, the curvature was kept constant. The local stress prediction gets etter when the local stresses are compared at the center of the plate. There are less oundary effects, and the curvature can e assumed constant for a unit cell. For investigation of the penalty of using the reduced stiffness matrix in the FSDT method for an unsymmetric ITPS panel, an unsymmetric panel with the same dimensions as mentioned aove was modeled and analyzed. The difference in this new model was that the top facesheet and wes were composed of Inconel 1, and the ottom facesheet was composed of an aluminum alloy (E :19 GPa, 1 :3). The local stresses were extracted at x a= and compared with the analytical stresses that resulted from the FSDT method y use of the reduced stiffness matrix. The comparison etween the FE results and the stress results otained from the FSDT method are illustrated in Figs. and 3. The results indicated that there was a % difference etween the analytical and FE results in the top facesheet when using the reduced stiffness matrix. The use of the reduced stiffness matrix in the FSDT method did not have any effect on the ottom facesheet and we stresses when compared with the results. The difference was much more noticeale in the top facesheet ecause of the local effects that resulted from the uniform pressure load. IV. Conclusions A -D plate solution for determining out-of-plane displacement of an orthotropic thick plate was estalished and verified with for a uniform pressure load and uniform thermal edge moments. Additional local facesheet deformations and stress were superimposed with the -D plate results to capture the local effects of the distriuted pressure load on the facesheets. Kirchhoff hypotheses and Euler Bernoulli plate equations were used for the local displacement and stress analysis of the facesheets and wes. The symmetric and unsymmetric plates resulted in a less than % prediction error in the plate s out-of-plane displacements when compared with the. The plate solution provided less than % prediction error results for ITPS panels with L=h ratios greater than 18. An ITPS panel with a low L=h introduced local oundary effects that contriuted to the panel s out-of-plane displacement. A 1-D eam analysis was done on the top facesheet to account for the local ending of the top facesheet due to its thin thickness. The local oundary effects that introduced local displacements and stresses were not fully captured with either the FSDT method or the CLPT method. Appendix A: Plate Equilirium Equations and Constitutive Relations The plate equilirium equations x Q x P z The plate constitutive relations are ) ( Mx M y M xy Qy D 11 D 1 D 1 D D Q xy Q y k C C 3( x;x y;y x;y y;x y w ;y x w ;x ) (A1) (A) (A3) Appendix B: Moments in the Half-Unit Cell In this Appendix, we provide the expressions for ending moments in various parts of the unit cell due to unit transverse force [see Eq. (1)]. The detailed derivations can e found in []: M AB yf y y f (B1) M BC ypy y p f (B) M DE y1 Fy P Rp Jd y p f (B3) M EG yry y f (B) M BE yff y cos Ppf y cos Hy sin (B) y s Appendix C: Local Bending Stress in Facesheets due to a Uniform Pressure Loading The moments in the equations shown elow are the result of the uniform pressure loading acting on the top facesheet. The moments were otained from a 1-D FE analysis. The moments were multiplied
11 MARTINEZ ET AL. 39 with the ending stiffness matrix of the facesheet to otain facesheet stress: M C EI L 3 L B L v B (C1) k i i 13 k i 1 ki 3 ki 11 i k i ki 11 D i i ki 1 ki 1 k i 11 D 11 i D C (D) (D8) M A EI L 3 L 1 v A L 1 B 1 (C) k i D i D C (D9) The ending moment distriution in the top facesheet consists of two equations: k i 1 ki 1 D 1 D i (D1) M BC y ym C P o y y p f (C3) k i 3 C (D11) M AB y yf B y M C P o L y y f (C) k i 13 C i (D1) Appendix D: Edge Moments In this section, the assumed solutions to the plate deflection with edge moments are provided. A relationship is also made to reduce the numer of unknown constants from 18 to. Ordinary differential equations to a -D plate with edge moments: D 11 B nxd C B n xd 1 D Cnx C A nx D CnxD C C n xd 1 D B nx C A n x C A nxc A n xc B nxc C n x A n x XN i1 B n x XN i1 C n x XN i1 i e ix i e ix i e ix (D1a) (D1) (D1c) (Da) (D) (Dc) A n x 1 e 1x e x 3 e 3x e x e x e x B n x 1 e 1x e x 3 e 3x e x e x e x C n x 1 e 1x e x 3 e 3x e x e x e x k i i 1 k i 3 ki ki 13 i k i ki 11 C i i ki 1 ki 1 (D3) (D) (D) (D) Acknowledgments This research is sponsored y a NASA grant under the Constellation University Institutes Project. The program manager is Claudia Mayer at NASA John H. Glenn Research Center at Lewis Field. References [1] Zhu, H., Design of Metallic Foams as Insulation in Thermal Protection Systems, Ph.D. Dissertation, Univ. of Florida, Gainesville, FL,. [] Blosser, M. L., Advanced Metallic Thermal Protection Systems for Reusale Launch Vehicles, Ph.D. Dissertation, Univ. of Virginia, Charlottesville, VA,. [3] Martinez, O., Sankar, B. V., Haftka, R., Blosser, M., and Bapanapalli, S. K., Micromechanical Analysis of a Composite Corrugated-Core Sandwich Panel for Integral Thermal Protection Systems, AIAA Journal, Vol., No. 9,, pp doi:1.1/1.9 [] Martinez, O., Sharma, A., Sankar, B. V., Haftka, R., and Blosser, M. L., Thermal Response Analysis of an Integrated Thermal Protection System for Future Space Vehicles, AIAA Journal, Vol. 8, No. 1, 1, pp doi:1.1/1.8 [] Fung, T. C., Tan, K. H., and Lok, T. S., Analysis of C-Core Sandwich Plate Decking, Proceedings of the 3rd International Offshore and Polar Engineering Conference, Mountain View, CA, Vol., 1993, pp. 9. [] Fung, T. C., Tan, K. H., and Lok, T. S., Elastic Constants for Z-Core Sandwich Panels, Journal of Structural Engineering, Vol. 1, No. 1, 199, pp doi:1.11/(asce)33-9(199)1:1(3) [] Liove, C., and Huka, R. E., Elastic Constants for Corrugated Core Sandwich Plates, NACA TN-89, 191. [8] Lok, T. S., Cheng, Q., and Heng, L., Equivalent Stiffness Parameters of Truss-Core Sandwich Panel, Proceedings of the Ninth International Offshore and Polar Engineering Conference, Vol., International Soc. of Offshore and Polar Engineers, Brest, France, 1999, pp [9] Valdevit, L., Hutchinson, J. W., and Evans, A. G., Structurally Optimized Sandwich Panels with Prismatic Cores, International Journal of Solids and Structures, Vol. 1, May, pp doi:1.11/j.ijsolstr... [1] Lok, T. S., and Cheng, Q., Elastic Stiffness Properties and Behavior of Truss-Core Sandwich Panel, Journal of Structural Engineering, Vol. 1, No., May, pp. 9. doi:1.11/(asce)33-9()1:() [11] Fung, T. C., Tan, K. H., and Lok, T. S., Shear Stiffness DQy for C-Core Sandwich Panels, Journal of Structural Engineering, Vol. 1, No. 8, 199, pp doi:1.11/(asce)33-9(199)1:8(98) [1] Liove, C., and Batdorf, S. B., A General Small Deflection Theory for Flat Sandwich Plates, NACA TN-1, 198. [13] Whitney, M., Structural Analysis of Laminated Anisotropic Plates, Technomic, Lancaster, PA, 198, Chaps. 9.
12 398 MARTINEZ ET AL. [1] Reddy, J. N., Mechanics of Laminated Composite Plates Theory and Analysis, CRC Press, Boca Raton, 199, Chaps.. [1] Martinez, O., Bapanapalli, S., Sankar, B. V., Haftka, R., and Blosser, M., Micromechanical Analysis of a Composite Truss Core Sandwich Panel for Integral Thermal Protection Systems, th AIAA/ASME/ ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference, Newport, RI, AIAA Paper -18, May. [1] Sankar, B. V., and Kim, N. H., Introduction to Finite Element Analysis and Design, Wiley, NY, 9, Chaps. 3. [1] Bapanapalli, S. K., Martinez, O., Sankar, B. V., Haftka, R. T., and Blosser, M. L., Analysis and Design of Corrugated-Core Sandwich Panels for Thermal Protection Systems of Space Vehicles, th AIAA/ ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference, Newport, RI, AIAA Paper -19, May. [18] Martinez, O., Sankar, B. V., and Haftka, R. T., Thermal Response Analysis of an Integral Thermal Protection System for Future Space Vehicles, ASME International Mechanical Engineering Congress and Exposition, Chicago, IL, American Soc. of Mechanical Engineers Paper -1, Fairfield, NJ, Nov.. [19] Martinez, O., Micromechanical Analysis and Design of an Integrated Thermal Protection System for Future Space Vehicles, Ph.D. Dissertation, Univ. of Florida, Gainesville, FL,. [] Timoshenko, S., and Woinowsky-Kreiger, S., Theory of Plates and Shells, Timoshenko, McGraw Hill, New York, 199, pp R. Ohayon Associate Editor
Homogenization of Integrated Thermal Protection System with Rigid Insulation Bars
5st AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference8th 2-5 April 200, Orlando, Florida AIAA 200-2687 Homogenization of Integrated Thermal Protection System with Rigid
More informationBuckling Behavior of Long Symmetrically Laminated Plates Subjected to Shear and Linearly Varying Axial Edge Loads
NASA Technical Paper 3659 Buckling Behavior of Long Symmetrically Laminated Plates Sujected to Shear and Linearly Varying Axial Edge Loads Michael P. Nemeth Langley Research Center Hampton, Virginia National
More informationComposite Plates Under Concentrated Load on One Edge and Uniform Load on the Opposite Edge
Mechanics of Advanced Materials and Structures, 7:96, Copyright Taylor & Francis Group, LLC ISSN: 57-69 print / 57-65 online DOI:.8/5769955658 Composite Plates Under Concentrated Load on One Edge and Uniform
More informationDavid A. Pape Department of Engineering and Technology Central Michigan University Mt Pleasant, Michigan
Session: ENG 03-091 Deflection Solutions for Edge Stiffened Plates David A. Pape Department of Engineering and Technology Central Michigan University Mt Pleasant, Michigan david.pape@cmich.edu Angela J.
More informationEffect of Thermal Stresses on the Failure Criteria of Fiber Composites
Effect of Thermal Stresses on the Failure Criteria of Fiber Composites Martin Leong * Institute of Mechanical Engineering Aalborg University, Aalborg, Denmark Bhavani V. Sankar Department of Mechanical
More informationHomework 6: Energy methods, Implementing FEA.
EN75: Advanced Mechanics of Solids Homework 6: Energy methods, Implementing FEA. School of Engineering Brown University. The figure shows a eam with clamped ends sujected to a point force at its center.
More informationPresented By: EAS 6939 Aerospace Structural Composites
A Beam Theory for Laminated Composites and Application to Torsion Problems Dr. BhavaniV. Sankar Presented By: Sameer Luthra EAS 6939 Aerospace Structural Composites 1 Introduction Composite beams have
More informationInternational Journal of Advanced Engineering Technology E-ISSN
Research Article INTEGRATED FORCE METHOD FOR FIBER REINFORCED COMPOSITE PLATE BENDING PROBLEMS Doiphode G. S., Patodi S. C.* Address for Correspondence Assistant Professor, Applied Mechanics Department,
More informationExact Shape Functions for Timoshenko Beam Element
IOSR Journal of Computer Engineering (IOSR-JCE) e-iss: 78-66,p-ISS: 78-877, Volume 9, Issue, Ver. IV (May - June 7), PP - www.iosrjournals.org Exact Shape Functions for Timoshenko Beam Element Sri Tudjono,
More informationHIGHER-ORDER THEORIES
HIGHER-ORDER THEORIES THIRD-ORDER SHEAR DEFORMATION PLATE THEORY LAYERWISE LAMINATE THEORY J.N. Reddy 1 Third-Order Shear Deformation Plate Theory Assumed Displacement Field µ u(x y z t) u 0 (x y t) +
More informationBending of Simply Supported Isotropic and Composite Laminate Plates
Bending of Simply Supported Isotropic and Composite Laminate Plates Ernesto Gutierrez-Miravete 1 Isotropic Plates Consider simply a supported rectangular plate of isotropic material (length a, width b,
More informationA HIGHER-ORDER BEAM THEORY FOR COMPOSITE BOX BEAMS
A HIGHER-ORDER BEAM THEORY FOR COMPOSITE BOX BEAMS A. Kroker, W. Becker TU Darmstadt, Department of Mechanical Engineering, Chair of Structural Mechanics Hochschulstr. 1, D-64289 Darmstadt, Germany kroker@mechanik.tu-darmstadt.de,
More informationNomenclature. Length of the panel between the supports. Width of the panel between the supports/ width of the beam
omenclature a b c f h Length of the panel between the supports Width of the panel between the supports/ width of the beam Sandwich beam/ panel core thickness Thickness of the panel face sheet Sandwich
More informationResearch Article Analysis Bending Solutions of Clamped Rectangular Thick Plate
Hindawi Mathematical Prolems in Engineering Volume 27, Article ID 7539276, 6 pages https://doi.org/.55/27/7539276 Research Article Analysis Bending Solutions of lamped Rectangular Thick Plate Yang Zhong
More informationParametric Investigation of Gas Permeability in Cross-Ply Laminates Using Finite Elements
AIAA JOURNAL Vol. 45, No. 4, April 7 Parametric Investigation of Gas Permeability in Cross-Ply Laminates Using Finite Elements Jianlong Xu and Bhavani V. Sankar University of Florida, Gainesville, Florida
More informationError Estimation and Error Reduction in Separable Monte Carlo Method
AIAA JOURNAL Vol. 48, No. 11, November 2010 Error Estimation and Error Reduction in Separable Monte Carlo Method Bharani Ravishankar, Benjamin P. Smarslok, Raphael T. Haftka, and Bhavani V. Sankar University
More informationAnalytical Strip Method for Thin Isotropic Cylindrical Shells
IOSR Journal of Mechanical and Civil Engineering (IOSR-JMCE) e-issn: 2278-1684,p-ISSN: 2320-334X, Volume 14, Issue 4 Ver. III (Jul. Aug. 2017), PP 24-38 www.iosrjournals.org Analytical Strip Method for
More informationFLEXURAL RESPONSE OF FIBER RENFORCED PLASTIC DECKS USING HIGHER-ORDER SHEAR DEFORMABLE PLATE THEORY
Asia-Pacific Conference on FRP in Structures (APFIS 2007) S.T. Smith (ed) 2007 International Institute for FRP in Construction FLEXURAL RESPONSE OF FIBER RENFORCED PLASTIC DECKS USING HIGHER-ORDER SHEAR
More informationACCURATE FREE VIBRATION ANALYSIS OF POINT SUPPORTED MINDLIN PLATES BY THE SUPERPOSITION METHOD
Journal of Sound and Vibration (1999) 219(2), 265 277 Article No. jsvi.1998.1874, available online at http://www.idealibrary.com.on ACCURATE FREE VIBRATION ANALYSIS OF POINT SUPPORTED MINDLIN PLATES BY
More information46th AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference April 2005 Austin, Texas
th AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics & Materials Conference - April, Austin, Texas AIAA - AIAA - Bi-stable Cylindrical Space Frames H Ye and S Pellegrino University of Cambridge, Cambridge,
More informationOpen Access Prediction on Deflection of V-core Sandwich Panels in Weak Direction
Send Orders for Reprints to reprints@benthamscience.net The Open Ocean Engineering Journal, 2013, 6, Suppl-1, M5) 73-81 73 Open Access Prediction on Deflection of V-core Sandwich Panels in Weak Direction
More informationStructural Dynamics Lecture Eleven: Dynamic Response of MDOF Systems: (Chapter 11) By: H. Ahmadian
Structural Dynamics Lecture Eleven: Dynamic Response of MDOF Systems: (Chapter 11) By: H. Ahmadian ahmadian@iust.ac.ir Dynamic Response of MDOF Systems: Mode-Superposition Method Mode-Superposition Method:
More informationHIGHER-ORDER THEORIES
HIGHER-ORDER THEORIES Third-order Shear Deformation Plate Theory Displacement and strain fields Equations of motion Navier s solution for bending Layerwise Laminate Theory Interlaminar stress and strain
More informationANALYTICAL AND EXPERIMENTAL ELASTIC BEHAVIOR OF TWILL WOVEN LAMINATE
ANALYTICAL AND EXPERIMENTAL ELASTIC BEHAVIOR OF TWILL WOVEN LAMINATE Pramod Chaphalkar and Ajit D. Kelkar Center for Composite Materials Research, Department of Mechanical Engineering North Carolina A
More informationUnit 15 Shearing and Torsion (and Bending) of Shell Beams
Unit 15 Shearing and Torsion (and Bending) of Shell Beams Readings: Rivello Ch. 9, section 8.7 (again), section 7.6 T & G 126, 127 Paul A. Lagace, Ph.D. Professor of Aeronautics & Astronautics and Engineering
More informationReview of Strain Energy Methods and Introduction to Stiffness Matrix Methods of Structural Analysis
uke University epartment of Civil and Environmental Engineering CEE 42L. Matrix Structural Analysis Henri P. Gavin Fall, 22 Review of Strain Energy Methods and Introduction to Stiffness Matrix Methods
More informationOnline publication date: 23 October 2010
This article was downloaded by: [informa internal users] On: 1 November 2010 Access details: Access Details: [subscription number 755239602] Publisher Taylor & Francis Informa Ltd Registered in England
More informationChapter 12 Plate Bending Elements. Chapter 12 Plate Bending Elements
CIVL 7/8117 Chapter 12 - Plate Bending Elements 1/34 Chapter 12 Plate Bending Elements Learning Objectives To introduce basic concepts of plate bending. To derive a common plate bending element stiffness
More informationBending, Vibration and Vibro-Acoustic Analysis of Composite Sandwich Plates with Corrugated Core. Rajesh Kumar Boorle
Bending, Vibration and Vibro-Acoustic Analysis of Composite Sandwich Plates with Corrugated Core by Rajesh Kumar Boorle A dissertation submitted in partial fulfillment of the requirements for the degree
More informationFinite element analysis of FRP pipelines dynamic stability
Boundary Elements and Other esh Reduction ethods XXXVIII 39 Finite element analysis of FRP pipelines dynamic stability D. G. Pavlou & A. I. Nergaard Department of echanical and Structural Engineering and
More informationLecture 8. Stress Strain in Multi-dimension
Lecture 8. Stress Strain in Multi-dimension Module. General Field Equations General Field Equations [] Equilibrium Equations in Elastic bodies xx x y z yx zx f x 0, etc [2] Kinematics xx u x x,etc. [3]
More informationAn Evaluation and Comparison of Models for Maximum Deflection of Stiffened Plates Using Finite Element Analysis
Marine Technology, Vol. 44, No. 4, October 2007, pp. 212 225 An Evaluation and Comparison of Models for Maximum Deflection of Stiffened Plates Using Finite Element Analysis Lior Banai 1 and Omri Pedatzur
More informationCHAPTER THREE SYMMETRIC BENDING OF CIRCLE PLATES
CHAPTER THREE SYMMETRIC BENDING OF CIRCLE PLATES * Governing equations in beam and plate bending ** Solution by superposition 1.1 From Beam Bending to Plate Bending 1.2 Governing Equations For Symmetric
More informationSIMILARITY METHODS IN ELASTO-PLASTIC BEAM BENDING
Similarity methods in elasto-plastic eam ending XIII International Conference on Computational Plasticity Fundamentals and Applications COMPLAS XIII E Oñate, DRJ Owen, D Peric and M Chiumenti (Eds) SIMILARIT
More informationKINGS COLLEGE OF ENGINEERING DEPARTMENT OF MECHANICAL ENGINEERING QUESTION BANK. Subject code/name: ME2254/STRENGTH OF MATERIALS Year/Sem:II / IV
KINGS COLLEGE OF ENGINEERING DEPARTMENT OF MECHANICAL ENGINEERING QUESTION BANK Subject code/name: ME2254/STRENGTH OF MATERIALS Year/Sem:II / IV UNIT I STRESS, STRAIN DEFORMATION OF SOLIDS PART A (2 MARKS)
More informationUNIT- I Thin plate theory, Structural Instability:
UNIT- I Thin plate theory, Structural Instability: Analysis of thin rectangular plates subject to bending, twisting, distributed transverse load, combined bending and in-plane loading Thin plates having
More informationDesign and Analysis of Metallic Thermal Protection System (MTPS)
International Journal of Scientific and Research Publications, Volume 3, Issue 2, February 2013 1 Design and Analysis of Metallic Thermal Protection System (MTPS) Suneeth Sukumaran*, Dr.S.H.Anilkumar**
More informationTABLE OF CONTENTS. Mechanics of Composite Materials, Second Edition Autar K Kaw University of South Florida, Tampa, USA
Mechanics of Composite Materials, Second Edition Autar K Kaw University of South Florida, Tampa, USA TABLE OF CONTENTS 1. INTRODUCTION TO COMPOSITE MATERIALS 1.1 Introduction... 1.2 Classification... 1.2.1
More informationDesign Variable Concepts 19 Mar 09 Lab 7 Lecture Notes
Design Variale Concepts 19 Mar 09 La 7 Lecture Notes Nomenclature W total weight (= W wing + W fuse + W pay ) reference area (wing area) wing aspect ratio c r root wing chord c t tip wing chord λ taper
More informationME 323 Examination #2
ME 33 Eamination # SOUTION Novemer 14, 17 ROEM NO. 1 3 points ma. The cantilever eam D of the ending stiffness is sujected to a concentrated moment M at C. The eam is also supported y a roller at. Using
More informationCOMPOSITE PLATE THEORIES
CHAPTER2 COMPOSITE PLATE THEORIES 2.1 GENERAL Analysis of composite plates is usually done based on one of the following the ries. 1. Equivalent single-layer theories a. Classical laminate theory b. Shear
More informationAN IMPROVED EFFECTIVE WIDTH METHOD BASED ON THE THEORY OF PLASTICITY
Advanced Steel Construction Vol. 6, No., pp. 55-547 () 55 AN IMPROVED EFFECTIVE WIDTH METHOD BASED ON THE THEORY OF PLASTICITY Thomas Hansen *, Jesper Gath and M.P. Nielsen ALECTIA A/S, Teknikeryen 34,
More informationChapter 4 Deflection and Stiffness
Chapter 4 Deflection and Stiffness Asst. Prof. Dr. Supakit Rooppakhun Chapter Outline Deflection and Stiffness 4-1 Spring Rates 4-2 Tension, Compression, and Torsion 4-3 Deflection Due to Bending 4-4 Beam
More informationDue Tuesday, September 21 st, 12:00 midnight
Due Tuesday, September 21 st, 12:00 midnight The first problem discusses a plane truss with inclined supports. You will need to modify the MatLab software from homework 1. The next 4 problems consider
More informationINTERNATIONAL JOURNAL OF APPLIED ENGINEERING RESEARCH, DINDIGUL Volume 2, No 2, 2011
Volume, No, 11 Copyright 1 All rights reserved Integrated Pulishing Association REVIEW ARTICLE ISSN 976 459 Analysis of free virations of VISCO elastic square plate of variale thickness with temperature
More informationLarge Amplitude Flexural Vibration of the Orthotropic Composite Plate Embedded with Shape Memory Alloy Fibers
Chinese Journal of Aeronautics 0(007) 5- Chinese Journal of Aeronautics www.elsevier.com/locate/cja Large Amplitude Flexural Viration of the Orthotropic Composite Plate Emedded with Shape Memory Alloy
More informationCHAPTER 14 BUCKLING ANALYSIS OF 1D AND 2D STRUCTURES
CHAPTER 14 BUCKLING ANALYSIS OF 1D AND 2D STRUCTURES 14.1 GENERAL REMARKS In structures where dominant loading is usually static, the most common cause of the collapse is a buckling failure. Buckling may
More informationComparison of Ply-wise Stress-Strain results for graphite/epoxy laminated plate subjected to in-plane normal loads using CLT and ANSYS ACP PrepPost
Comparison of Ply-wise Stress-Strain results for graphite/epoxy laminated plate subjected to in-plane normal loads using CLT and ANSYS ACP PrepPost 1 Mihir A. Mehta, 2 Satyen D. Ramani 1 PG Student, Department
More informationMODELING THE OUT-OF-PLANE BENDING BEHAVIOR OF RETROFITTED URM WALLS. Abstract. Introduction
MOELING THE OUT-OF-PLANE ENING EHAVIOR OF RETROFITTE URM WALLS J. I. Velázuez imas, Professor, Facultad de Ingeniería, Universidad Autonoma de Sinaloa, Mexico M.R. Ehsani, Professor CEEM epartment, The
More informationBUCKLING COEFFICIENTS FOR SIMPLY SUPPORTED, FLAT, RECTANGULAR SANDWICH PANELS UNDER BIAXIAL COMPRESSION
U. S. FOREST SERVICE RESEARCH PAPER FPL 135 APRIL 1970 BUCKLING COEFFICIENTS FOR SIMPLY SUPPORTED, FLAT, RECTANGULAR SANDWICH PANELS UNDER BIAXIAL COMPRESSION FOREST PRODUCTS LABORATORY, FOREST SERVICE
More informationComposite Structural Mechanics using MATLAB
Session 2520 Composite Structural Mechanics using MATLAB Oscar Barton, Jr., Jacob B. Wallace United States Naval Academy Annapolis, Md 21402 Abstract In this paper MATLAB is adopted as the programming
More informationAccepted Manuscript. R.C. Batra, J. Xiao S (12) Reference: COST Composite Structures. To appear in:
Accepted Manuscript Finite deformations of curved laminated St. Venant-Kirchhoff beam using layerwise third order shear and normal deformable beam theory (TSNDT) R.C. Batra, J. Xiao PII: S0263-8223(12)00486-2
More informationShafts. Fig.(4.1) Dr. Salah Gasim Ahmed YIC 1
Shafts. Power transmission shafting Continuous mechanical power is usually transmitted along and etween rotating shafts. The transfer etween shafts is accomplished y gears, elts, chains or other similar
More informationSection 8.5. z(t) = be ix(t). (8.5.1) Figure A pendulum. ż = ibẋe ix (8.5.2) (8.5.3) = ( bẋ 2 cos(x) bẍ sin(x)) + i( bẋ 2 sin(x) + bẍ cos(x)).
Difference Equations to Differential Equations Section 8.5 Applications: Pendulums Mass-Spring Systems In this section we will investigate two applications of our work in Section 8.4. First, we will consider
More informationMIXED RECTANGULAR FINITE ELEMENTS FOR PLATE BENDING
144 MIXED RECTANGULAR FINITE ELEMENTS FOR PLATE BENDING J. N. Reddy* and Chen-Shyh-Tsay School of Aerospace, Mechanical and Nuclear Engineering, University of Oklahoma, Norman, Oklahoma The paper describes
More informationModule 5: Laminate Theory Lecture 19: Hygro -thermal Laminate Theory. Introduction: The Lecture Contains: Laminate Theory with Thermal Effects
Module 5: Laminate Theory Lecture 19: Hygro -thermal Laminate Theory Introduction: In this lecture we are going to develop the laminate theory with thermal and hygral effects. Then we will develop the
More informationStress Analysis on Bulkhead Model for BWB Heavy Lifter Passenger Aircraft
The International Journal Of Engineering And Science (IJES) Volume 3 Issue 01 Pages 38-43 014 ISSN (e): 319 1813 ISSN (p): 319 1805 Stress Analysis on Bulkhead Model for BWB Heavy Lifter Passenger Aircraft
More informationFlexural-Torsional Buckling of General Cold-Formed Steel Columns with Unequal Unbraced Lengths
Proceedings of the Annual Stability Conference Structural Stability Research Council San Antonio, Texas, March 21-24, 2017 Flexural-Torsional Buckling of General Cold-Formed Steel Columns with Unequal
More informationGeneric Strategies to Implement Material Grading in Finite Element Methods for Isotropic and Anisotropic Materials
University of Nebraska - Lincoln DigitalCommons@University of Nebraska - Lincoln Engineering Mechanics Dissertations & Theses Mechanical & Materials Engineering, Department of Winter 12-9-2011 Generic
More informationP = ρ{ g a } + µ 2 V II. FLUID STATICS
II. FLUID STATICS From a force analysis on a triangular fluid element at rest, the following three concepts are easily developed: For a continuous, hydrostatic, shear free fluid: 1. Pressure is constant
More informationAssessment of the Buckling Behavior of Square Composite Plates with Circular Cutout Subjected to In-Plane Shear
Assessment of the Buckling Behavior of Square Composite Plates with Circular Cutout Sujecte to In-Plane Shear Husam Al Qalan 1)*, Hasan Katkhua 1) an Hazim Dwairi 1) 1) Assistant Professor, Civil Engineering
More informationExact Free Vibration of Webs Moving Axially at High Speed
Eact Free Viration of Wes Moving Aially at High Speed S. HATAMI *, M. AZHARI, MM. SAADATPOUR, P. MEMARZADEH *Department of Engineering, Yasouj University, Yasouj Department of Civil Engineering, Isfahan
More informationCHAPTER 5 PROPOSED WARPING CONSTANT
122 CHAPTER 5 PROPOSED WARPING CONSTANT 5.1 INTRODUCTION Generally, lateral torsional buckling is a major design aspect of flexure members composed of thin-walled sections. When a thin walled section is
More informationStatic & Free Vibration Analysis of an Isotropic and Orthotropic Thin Plate using Finite Element Analysis (FEA)
International Journal of Current Engineering and Technology E-ISSN 2277 4106, P-ISSN 2347 5161 2015 INPRESSCO, All Rights Reserved Available at http://inpressco.com/category/ijcet Research Article Static
More information4. BEAMS: CURVED, COMPOSITE, UNSYMMETRICAL
4. BEMS: CURVED, COMPOSITE, UNSYMMETRICL Discussions of beams in bending are usually limited to beams with at least one longitudinal plane of symmetry with the load applied in the plane of symmetry or
More informationGerald Allen Cohen, 83, passed away Oct. 1, 2014, at his home in Laguna Beach.
Dr Gerald Allen Cohen (1931-2014) Ring-stiffened shallow conical shell designed with the use of FASOR for NASA s Viking project in the 1970s. (from NASA TN D-7853, 1975, by Walter L. Heard, Jr., Melvin
More informationOptimum Height of Plate Stiffener under Pressure Effect
The st Regional Conference of Eng. Sci. NUCEJ Spatial ISSUE vol., No.3, 8 pp 459-468 Optimum Height of Plate Stiffener under Pressure Effect Mazin Victor Yousif M.Sc Production Engineering University of
More informationFinite Element Method in Geotechnical Engineering
Finite Element Method in Geotechnical Engineering Short Course on + Dynamics Boulder, Colorado January 5-8, 2004 Stein Sture Professor of Civil Engineering University of Colorado at Boulder Contents Steps
More informationAPPLICATION OF THE GALERKIN-VLASOV METHOD TO THE FLEXURAL ANALYSIS OF SIMPLY SUPPORTED RECTANGULAR KIRCHHOFF PLATES UNDER UNIFORM LOADS
Nigerian Journal of Technology (NIJOTECH) Vol. 35, No. 4, October 2016, pp. 732 738 Copyright Faculty of Engineering, University of Nigeria, Nsukka, Print ISSN: 0331-8443, Electronic ISSN: 2467-8821 www.nijotech.com
More informationUnit 18 Other Issues In Buckling/Structural Instability
Unit 18 Other Issues In Buckling/Structural Instability Readings: Rivello Timoshenko Jones 14.3, 14.5, 14.6, 14.7 (read these at least, others at your leisure ) Ch. 15, Ch. 16 Theory of Elastic Stability
More informationExtensional and Flexural Waves in a Thin-Walled Graphite/Epoxy Tube * William H. Prosser NASA Langley Research Center Hampton, VA 23665
Extensional and Flexural Waves in a Thin-Walled Graphite/Epoxy Tube * William H. Prosser NASA Langley Research Center Hampton, VA 23665 Michael R. Gorman Aeronautics and Astronautics Naval Postgraduate
More information2766. Differential quadrature method (DQM) for studying initial imperfection effects and pre- and post-buckling vibration of plates
2766. Differential quadrature method (DQM) for studying initial imperfection effects and pre- and post-buckling vibration of plates Hesam Makvandi 1, Shapour Moradi 2, Davood Poorveis 3, Kourosh Heidari
More informationFinite Element Analysis Prof. Dr. B. N. Rao Department of Civil Engineering Indian Institute of Technology, Madras. Module - 01 Lecture - 13
Finite Element Analysis Prof. Dr. B. N. Rao Department of Civil Engineering Indian Institute of Technology, Madras (Refer Slide Time: 00:25) Module - 01 Lecture - 13 In the last class, we have seen how
More informationFig. 1. Circular fiber and interphase between the fiber and the matrix.
Finite element unit cell model based on ABAQUS for fiber reinforced composites Tian Tang Composites Manufacturing & Simulation Center, Purdue University West Lafayette, IN 47906 1. Problem Statement In
More informationComposites Design and Analysis. Stress Strain Relationship
Composites Design and Analysis Stress Strain Relationship Composite design and analysis Laminate Theory Manufacturing Methods Materials Composite Materials Design / Analysis Engineer Design Guidelines
More informationAircraft Structures Beams Torsion & Section Idealization
Universit of Liège Aerospace & Mechanical Engineering Aircraft Structures Beams Torsion & Section Idealiation Ludovic Noels omputational & Multiscale Mechanics of Materials M3 http://www.ltas-cm3.ulg.ac.be/
More informationRiveted Joints and Linear Buckling in the Steel Load-bearing Structure
American Journal of Mechanical Engineering, 017, Vol. 5, No. 6, 39-333 Availale online at http://pus.sciepu.com/ajme/5/6/0 Science and Education Pulishing DOI:10.1691/ajme-5-6-0 Riveted Joints and Linear
More informationFE Modeling of Fiber Reinforced Polymer Structures
WCCM V Fifth World Congress on Computational Mechanics July 7 12, 2002, Vienna, Austria Eds.: H.A. Mang, F.G. Rammerstorfer, J. Eerhardsteiner FE Modeling of Fier Reinforced Polymer Structures W. Wagner
More informationEsben Byskov. Elementary Continuum. Mechanics for Everyone. With Applications to Structural Mechanics. Springer
Esben Byskov Elementary Continuum Mechanics for Everyone With Applications to Structural Mechanics Springer Contents Preface v Contents ix Introduction What Is Continuum Mechanics? "I Need Continuum Mechanics
More informationOverview of Probabilistic Modeling of Woven Ceramic Matrix Composites ABSTRACT
ABSTRACT An overview of current research efforts in understanding the cause of variability in the thermo-mechanical properties of woven ceramic matrix composites is presented. Statistical data describing
More informationParametric study on the transverse and longitudinal moments of trough type folded plate roofs using ANSYS
American Journal of Engineering Research (AJER) e-issn : 2320-0847 p-issn : 2320-0936 Volume-4 pp-22-28 www.ajer.org Research Paper Open Access Parametric study on the transverse and longitudinal moments
More informationNONLINEAR STRUCTURAL DYNAMICS USING FE METHODS
NONLINEAR STRUCTURAL DYNAMICS USING FE METHODS Nonlinear Structural Dynamics Using FE Methods emphasizes fundamental mechanics principles and outlines a modern approach to understanding structural dynamics.
More informationUnit 13 Review of Simple Beam Theory
MIT - 16.0 Fall, 00 Unit 13 Review of Simple Beam Theory Readings: Review Unified Engineering notes on Beam Theory BMP 3.8, 3.9, 3.10 T & G 10-15 Paul A. Lagace, Ph.D. Professor of Aeronautics & Astronautics
More informationFLEXIBILITY METHOD FOR INDETERMINATE FRAMES
UNIT - I FLEXIBILITY METHOD FOR INDETERMINATE FRAMES 1. What is meant by indeterminate structures? Structures that do not satisfy the conditions of equilibrium are called indeterminate structure. These
More informationEFFECTS OF THERMAL STRESSES AND BOUNDARY CONDITIONS ON THE RESPONSE OF A RECTANGULAR ELASTIC BODY MADE OF FGM
Proceedings of the International Conference on Mechanical Engineering 2007 (ICME2007) 29-31 December 2007, Dhaka, Bangladesh ICME2007-AM-76 EFFECTS OF THERMAL STRESSES AND BOUNDARY CONDITIONS ON THE RESPONSE
More informationDynamic and buckling analysis of FRP portal frames using a locking-free finite element
Fourth International Conference on FRP Composites in Civil Engineering (CICE8) 22-24July 8, Zurich, Switzerland Dynamic and buckling analysis of FRP portal frames using a locking-free finite element F.
More informationVIBRATION AND DAMPING ANALYSIS OF FIBER REINFORCED COMPOSITE MATERIAL CONICAL SHELLS
VIBRATION AND DAMPING ANALYSIS OF FIBER REINFORCED COMPOSITE MATERIAL CONICAL SHELLS Mechanical Engineering Department, Indian Institute of Technology, New Delhi 110 016, India (Received 22 January 1992,
More informationUNIVERSITY OF SASKATCHEWAN ME MECHANICS OF MATERIALS I FINAL EXAM DECEMBER 13, 2008 Professor A. Dolovich
UNIVERSITY OF SASKATCHEWAN ME 313.3 MECHANICS OF MATERIALS I FINAL EXAM DECEMBER 13, 2008 Professor A. Dolovich A CLOSED BOOK EXAMINATION TIME: 3 HOURS For Marker s Use Only LAST NAME (printed): FIRST
More informationMechanics of Materials II. Chapter III. A review of the fundamental formulation of stress, strain, and deflection
Mechanics of Materials II Chapter III A review of the fundamental formulation of stress, strain, and deflection Outline Introduction Assumtions and limitations Axial loading Torsion of circular shafts
More informationPassive Damping Characteristics of Carbon Epoxy Composite Plates
Journal of Materials Science and Engineering A 6 (-) 35-4 doi:.765/6-63/6.-.5 D DAVID PUBLISHING Passive Damping Characteristics of Carbon Epoxy Composite Plates Dileep Kumar K * and V V Subba Rao Faculty
More informationThin-walled Box Beam Bending Distortion Analytical Analysis
Journal of Mechanical Engineering Vol 14(1), 1-16, 217 Thin-walled Box Beam Bending Distortion Analytical Analysis A Halim Kadarman School of Aerospace Engineering Universiti Sains Malaysia, 143 Nibong
More informationSimulation of inelastic cyclic buckling behavior of steel box sections
Computers and Structures 85 (27) 446 457 www.elsevier.com/locate/compstruc Simulation of inelastic cyclic uckling ehavior of steel ox sections Murat Dicleli a, *, Anshu Mehta a Department of Engineering
More informationME 354 MECHANICS OF MATERIALS LABORATORY STRESSES IN STRAIGHT AND CURVED BEAMS
ME 354 MECHNICS OF MTERILS LBORTORY STRESSES IN STRIGHT ND CURVED BEMS OBJECTIVES January 2007 NJS The ojectives of this laoratory exercise are to introduce an experimental stress analysis technique known
More informationEffect of magnetostrictive material layer on the stress and deformation behaviour of laminated structure
IOP Conference Series: Materials Science and Engineering PAPER OPEN ACCESS Effect of magnetostrictive material layer on the stress and deformation behaviour of laminated structure To cite this article:
More informationBI-STABLE COMPOSITE SHELLS
AIAA 2-1385 BI-STABLE COMPOSITE SHELLS K. Iqbal and S. Pellegrino Department of Engineering, University of Cambridge, Trumpington Street, Cambridge, CB2 1PZ, U.K. Abstract This paper is concerned with
More informationLaminated Beams of Isotropic or Orthotropic Materials Subjected to Temperature Change
United States Department of Agriculture Forest Service Forest Products Laboratory Research Paper FPL 375 June 1980 Laminated Beams of Isotropic or Orthotropic Materials Subjected to Temperature Change
More informationBending Analysis of Symmetrically Laminated Plates
Leonardo Journal of Sciences ISSN 1583-0233 Issue 16, January-June 2010 p. 105-116 Bending Analysis of Symmetrically Laminated Plates Bouazza MOKHTAR 1, Hammadi FODIL 2 and Khadir MOSTAPHA 2 1 Department
More informationCIV-E1060 Engineering Computation and Simulation Examination, December 12, 2017 / Niiranen
CIV-E16 Engineering Computation and Simulation Examination, December 12, 217 / Niiranen This examination consists of 3 problems rated by the standard scale 1...6. Problem 1 Let us consider a long and tall
More informationA New Method of Analysis of Continuous Skew Girder Bridges
A New Method of Analysis of Continuous Skew Girder Bridges Dr. Salem Awad Ramoda Hadhramout University of Science & Technology Mukalla P.O.Box : 50511 Republic of Yemen ABSTRACT One three girder two span
More informationAEROELASTIC ANALYSIS OF SPHERICAL SHELLS
11th World Congress on Computational Mechanics (WCCM XI) 5th European Conference on Computational Mechanics (ECCM V) 6th European Conference on Computational Fluid Dynamics (ECFD VI) E. Oñate, J. Oliver
More information